se21 paper-cellular floors/long span beams in fire tests...

18 The Structural Engineer 89 (21) 1 November 2011

Synopsis

As part of an international major research initiative, dealing with thebehaviour of long span cellular beams in steel framed buildingsunder fire, a large-scale fire test incorporating 15m long cellularbeams was carried out. The test incorporated unprotectedsecondary cellular steel beams acting compositely with thesupporting floor slab. The floorplate in its entirety was designed tocarry the load with unprotected beams, when subjected to asevere fire, by utilising membrane action of the floor slab. Theoverall structure performed very well supporting the full appliedstatic load for the duration of the test. The unprotected cellularsteel beams were subjected to distortional buckling, with only thetop tee providing any support through catanary action. The testsupported the assumptions adopted in the structural design

approach and provided an accurate estimate of the strength of thefloorplate. Comparison of the recorded time-temperaturerelationship of the fire with the design method presented in theEurocodes shows that the code under-predicts the severity of thefire, although this was compensated to some extent by theconservative assumptions embedded within the structural model.

Introduction

To reduce fabrication costs and provide greater flexibility withinbuildings, horizontal members need to span greater distances andthe number of columns need to be kept to a minimum. In terms ofsteel-framed buildings greater efficiency, and thus longer spans,can be obtained with the steel beams acting compositely with thesupported floorplate. However, one disadvantage of longer spansis the need to increase the beam’s overall depth. If the services arethen hung in a zone below the beams the depth of the floor zoneincreases, which will result in the overall building getting higher withan increase in cladding costs. Alternatively if the overall buildingheight is restricted, due to possible planning constraints, the totalnumber of floors within the building will be reduced with an overallreduction in lettable floor space. One way to overcome thisproblem is to use cellular beams (CBs) with uniformly spacedholes, as shown in Fig 1(a), with ducts, pipes and other servicespassing through the openings in the web, thus reducing the heightof the overall floor zone. The use of CBs also has the advantage ofreducing the amount of steel material used and the overall deadweight of the steelwork.

There has previously been significant research work1–3

investigating the detailed and complex behaviour of CBs at normaltemperature, which has included understanding various failuremodes comprising web-post buckling, Vierendeel bending, web-weld failure, overall lateral-torsional buckling, distortional bucklingand flexural bending. The understanding of CBs in fire conditionshas also recently been enhanced, through modelling4–6 and anumber of fire tests7, to ensure adequate safety under fireconditions. This has also included tests investigating the actualdetailed performance of the fire protection when applied to CBs8.Although, without doubt, a greater understanding of theperformance of CBs under fire conditions has recently beenobtained this has been limited to test data and observations fromsmall-scale, and arguably unrealistic, fire tests.

Following the fire tests on a full sized steel-framed office buildingat Cardington in the UK9,10, a simple design method for beams and

Paper

Full-scale fire test on a composite floor slabincorporating long span cellular steel beams

A. Nadjai, BEng, MSc, PhD, MIFireEProfessor of Fire Structural Engineering, University of Ulster, UK

C. G. Bailey, BEng, PhD, FICE, MIStructE, MIFireEProfessor of Structural Engineering, University of Manchester, UK

Eur.-Ing. O. Vassart, PhDSenior Research Engineer ArcelorMittal, G.-D. of Luxemburg

S. Han, BEng, MSc, PhDResearch Associate, University of Ulster, UK

Eur.-Ing. B. Zhao, PhD Head of Fire Engineering R&D department, CTICM, France

Eur.-Ing. M. HawesTechnical Director of ASDWestok, UK

Eur.-Ing. J. -M. Franssen, PhDProfessor at the University of Liège, Belgium

I. Simms, BEng(Hons), PhD, MICESenior Engineer, Steel Construction Institute, UK

Keywords: Full scale tests, Fire tests, Cellular beams, Steel, Frames, Buildings, Floors,

Composite construction, Slabs

Received: 11/10: Modified: 11/10-6/11; Accepted: 07/11

© A. Nadjai, C. G. Bailey, O. Vassart, S. Han, B. Zhoa, M. Hawes, J. -M. Franssen & I. Simms

1(a) Test compartment with long unprotected cellular beams 1(b) Steel structural layout

SE21 paper-Cellular floors/Long span beams in fire tests_Layout 3 26/10/2011 16:47 8

The Structural Engineer 89 (21) 1 November 2011 19

floor systems in fire was developed by Bailey11-14. The Baileymethod allows a number of steel beams, within a given floorplate,to be left unprotected by indentifying only those beams thatactually require applied fire protection and utilising the inherentstrength of the system through membrane action within thecomposite floor system. The Bailey method is presented in designguides15,16 and software published by the Building ResearchEstablishment and the Steel Construction Institute.

Although the tests at Cardington provided a significant stepforward in terms of understanding the structural behaviour of steel-framed buildings under fire, the maximum span of the beams inthe tests was limited to 9m, whereas spans of 15m are morecommon when cellular beams are used. In addition, there hasbeen concern expressed that the localised modes of behaviourobserved in CBs, such as web-buckling, may influence the globalbehaviour of the floor slab and possibly inhibit the development ofmembrane action, which the design method relies on.

To further validate the Bailey method and enhance theunderstanding of the behaviour of CBs under fire conditions, theUniversity of Ulster, under the direction of Professor Nadjai,designed and constructed a large-scale natural fire test on acomposite floor slab supported by long-span cellular beams. Thedata from the test will provide further confidence in the use of theexisting design guidance, provide validation for the use ofadvanced models17 and highlight the inherent fire resistance of thesystem.

Test details

The tested floorplate was 9.6m by 15.6m supported on a steelframe spanning 9m by 15m between four corner columns (Fig 1b).The cellular beams were positioned on gridlines 1, 4, B, C and Das primary and secondary beams of the structure (Fig 1b). Thedimensions of the beams are shown in Figs 1b and 1c. Theunprotected secondary Beams 4 and 5 also had an elongated

web opening at the centre of their span (Figs 1a and 5). All theconnections between the secondary and primary beams andbetween the beams and columns were simple fin plates, exceptfor the connections between Beam 5 and the primary beams,which were partial end-plates (Fig 1d). The structure was designedat the time using BS 5950-118 and BS 5950-319.

The enclosed compartment was 9.2m by 15.6m, with aninternal floor-to-soffit height of 2.88m. The surrounding walls wereconstructed using 7N/mm2 blockwork, with three openings, each1.5m by 3m (Fig 3e), along gridline. The surrounding compartmentwalls along gridlines 1, 4 and D were not fixed to the compositefloor at the top which allowed free vertical movement of thefloorplate along these boundaries. The front façade, with openings,was constructed such that the wall was extended up to theunderside of the solid beam along gridline A, allowing no verticaldeflection of the beam along this gridline (Fig 1c). The frame wasbraced in the horizontal direction at the following locations;Column A1 was braced in both lateral directions, Column A4 wasbraced laterally parallel to gridline 4 and Column D1 was bracedlaterally parallel to gridline D. Bracing was provided using adiagonal CHS, as shown in Fig 1e.

All the columns, and the solid beam along gridline A, wereprotected using commercially available 20mm thick fire board witha standard fire resistance period of 2h. The perimeter CBs ongridlines 1, 4, and D were protected using a ceramic fibre (see Fig2), which also provided a standard fire resistance period of 2h. Thefire protection was fitted using an approved contractor, followingthe manufacturer’s specification. Plasterboard, 15mm thick, wasalso used to cover the inner face of the boundary walls to reduceheat loss through the blockwork (Fig 2).

The concrete composite slab was 120mm thick and compriseda 51mm deep, 1mm thick, Holorib steel deck (HR51/150), normal-weight concrete and mesh steel reinforcement. The dovetail steeldeck had a measured tensile strength of 327N/mm2. The welded

1(e) Elevation of the compartment façade

1(c) Detail information of the steel sections

1(d) Beam connections before protection applied to the primary beams

2 Fibre and plasterboard protection used inside the compartment

SE21 paper-Cellular floors/Long span beams in fire tests_Layout 3 26/10/2011 16:47 Page 19


wire A393 mesh reinforcement (Fig 3) comprised 10mm diameterribbed bars at 200mm centres, with nominal yield strength of500N/mm2, which was specified using the Bailey Method15, basedon the design parametric fire curve. The mesh reinforcement had aminimum lap length of 400mm and covered with 40mm thicknessof concrete. The concrete mix design (for 1m³) comprised: 320kgOPC, 918kg 10mm limestone, 691kg sharp sand, 380kg 6mmlimestone, 30kg grey (recycled) water and 142kg cold (tap) water.No additives or air entraining agent was used in the concretemixture. The measured average concrete compressive cubestrength was 50N/mm2 on the day of test.

The slab was exposed to an external environment prior totesting, with no protection from the Northern Ireland weather,resulting in a measured moisture content of 6.4% (by weight) at thetime of the test. It can be argued that at 6.4% the moisture contentis unrealistically high, which can result in spalling of the concrete ora reduction in the temperature of floor slab compared to astructure which has more realistic moisture content. In terms ofspalling the metal deck will alleviate, to some extent, thisdetrimental effect since it will hold the spalled concrete in place,provided that the deck does not significantly de-bond and a large

gap is not created between the deck and concrete. The highmoisture content will result in lower temperatures through theconcrete slab resulting in lower temperatures of the meshreinforcement. To consider the effect of the moisture content onthe temperatures and resulting structural behaviour, detailedcomputer modelling is currently being conducted by the authors.

Full interaction between the slab and beams was achieved usingshear connectors, of 19mm diameter and 95mm height, placed at200mm centres along the beams (using BS 5950-318). Therequirement for U-bar reinforcement around the slab’s perimeter(as shown in Fig 4) is not a special requirement for fire design, butwas needed to ensure correct reinforcement detailing for ambientdesign based on the recommendations given in the SteelConstruction Institute Publications P300 and AD32520, 21. The U-bars were 10mm diameter and placed with 30mm cover to theedge of the slab, as shown in Fig 4.

Instrumentation

Extensive instrumentation devices were placed throughout thecompartment to measure the atmosphere temperatures,temperature distribution through the composite floor, thetemperature of the protected and unprotected cellular beams, andthe vertical and horizontal displacements. The locations of themeasurements taken are shown in Figs 4 and 5. A free-standingsteel structure was built around the compartment to create a

DescriptionkN/m2

Characteristic Load(kN/m2)

Load Factor atFLS

Design Load atFLS

Partition 1.0 1.0 1.0

Services &Finishes 0.5 1.0 0.5

Live Load 3.5 0.5 1.75

Total 3.25

5 Thermocouple locations on unprotected Beam 4 (Gridline B)

6(a) Vertical static load

3 Mesh reinforcement and steel decking before concrete casting

4 Locations of measurement positions for deflections and temperatures throughout the slab

6(b) Wooden cribs used for the fire load

Table 1 Load table



reference outer frame, allowing the correct measurement of verticaland horizontal displacements (Figs 3 and 6a). A total of 350thermocouples were used to monitor the temperatures and a totalof 17 transducers were used to measure the variousdisplacements. The transducers were attached to a free-standingouter reference frame and were insulated, where required, toensure that and heat effects to the transducers were eliminated.

Design loads

The design load was based on a characteristic live load of3.5kN/m2 together with a partition load of 1.0kN/m2 and a servicesand finishes load of 0.5kN/m2. The partial load factors used for theFire Limit State (FLS) correspond to the values given in the UKNational Annex to BS EN199022 for office buildings. The resultingapplied load was 3.25kN/m2, as shown in Table 1 on page 20.

The applied load was achieved using 44 sandbags (eachweighing 1t) evenly positioned over the floorplate, as shown in Fig6a, providing a load of 3.25kN/m2. The self weight of the slab,which was 120mm thick, was calculated as 2.90kN/m2, creating atotal load of 6.15kN/m2.

Design of the fire

The natural fire was designed using the parametric time-temperature curves in Annex A of BSEN1991-1-223. The fire loadcomprised 45 standard (1m × 1m × 0.5m high) wooden cribs, builtusing 50mm x 50mm x 1000mm wooden battens, positionedevenly around the compartment (Fig 6b). The fire load wasequivalent to 40kg of wood per square metre of floor area.Assuming a calorific value of 17.5MJ/kg for wood, the fire loaddensity for the tested compartment was 587MJ/m2. The fire loadused was slightly higher than the office design fire load of570MJ/m2 (80% fractile) given in the UK National Annex24 and in

BS PD 7974-125. The adopted fire load was also significantlyhigher than the value of 511MJ/m2, which is the Europeanminimum recommended value given in Annex E of BS EN 1991-1-223. Each wooden crib was connected to its neighbour by a mildsteel channel section, which contained a porous fibre board.Approximately 30min before ignition, 20 litres of paraffin waspoured into the channels, to ensure rapid fire development withinthe compartment.

Based on the actual fire load, ventilation openings, compartmentgeometry and representative compartment boundary insulationmaterials (b factor taken as 1104J/m2s1/2K), the parametric firecurve can be defined using the method outlined in BS EN 1991-1-223. The resulting parametric curve is shown in Fig. 7 with apredicted maximum temperature of 921°C occurring at 52 min.The OZone V2 model26, which is a simple model that defines auniform zone temperature of the gases as a function of time, byconsidering the conservation of mass and energy in the firecompartment17, was also used to predict the temperature-timeresponse of the test compartment. The OZone V2 model gave amaximum temperature of 951°C at 62min, as shown in Fig 7.

Test observations

The fire was ignited at a single point in the middle of thecompartment (Fig 8). After 7 minutes it was decided to ignite thecribs at two further locations close to gridlines 1 and 4, and thenthe fire was left to develop by itself (Fig 9), without any interference.The maximum recorded atmosphere temperature of 1055°Coccurred in the centre of the compartment, 500mm below theceiling after 75min. Fig 7 shows the recorded average atmospheretime-temperature compared to the calculated design parametricfire curve and the OZone V2 model, using the test parameters. Ifthe OZone and parametric curves are ‘shifted’ for an assumed

9 Developed compartment fire 10 Deflection of the slab/unprotected beam following the fire

7 Design parametric fire curve, OZone model and measured atmosphere temperatures

8 Ignition of the fire



delay of 5 and 15min of ignition respectively, we can see that bothcurves are more aligned with the measured fire temperatures.However, it can be seen that both design curves areunconservative in terms of maximum temperature and the durationover which the maximum temperature occurs. The parametriccurve has been shown to be unconservative in previous full-scaletests27, and further work is needed to address this issue.

Beam/slab deflection

Under fire conditions the deflection of the unprotected, axiallyunrestrained, composite steel beams (Fig 10) predominatelycomprises two parts; thermal bowing and mechanical deflection.Deflection due to thermal bowing is caused by the non-uniformtemperature distribution through the steel beam and theconnected composite slab. The mechanical deflection is due tothe decrease in stiffness and strength of the structural material asthe temperatures increase. At low temperatures (less than 400°C),the beam deflection is predominantly due to thermal bowing. Athigher temperatures, mechanical deflection will dominate and thedeflection increases at a faster rate.

The maximum recorded steel temperature of 1053°C occurredafter 77min at the centre span of Beams 4 and 5 (Fig 11). Themaximum temperature occurred on the bottom flange below theelongated opening. Fig 12 shows the temperature distribution atthe critical part of the unprotected CBs. It is worth noting that thetemperatures are non-uniform across the web despite the beamsbeing unprotected and the long duration of the fire. Thetemperature of the top flange of the beams is lower, as expected,due to the heat sink effect of the supporting concrete slab. At amaximum temperature of 1053°C the steel has lost 97% of itsstrength and stiffness27 and is contributing little to the load bearingcapacity of the floor system.

With increasing temperatures on the unprotected CBs (Fig 11), itwas observed that post web buckling occurred initially. The

composite action between the CBs and slab prevented twisting ofthe beam as a whole. The tendency for the bottom flange todisplace laterally caused bending of the beam’s web leading tooverall distortional buckling, as shown in Fig 10. At this stage theunprotected steel temperatures were approximately 800°C andonly the top flange was considered to be providing support to theslab by acting as a catenary (Fig 10). The temperature of the meshreinforcement, above the beams reached a maximum of 375°C at95min, as shown in Fig 13a which was well into the cooling stagesof the fire. Fig 13b shows the maximum recorded temperature ofthe mesh reinforcement between the beams, where again themaximum temperature occurred during the cooling stages of thefire. The temperature in the concrete slab continues to rise afterthe maximum atmosphere temperature, which occurred at 75mins.The recorded temperatures of the shear studs are shown in Fig13b, where the maximum temperature reached 585°C. Althoughthe shear stud temperature is high the amount of horizontal shearrequired reduces as the unprotected beams increase intemperature and lose strength and stiffness. There was no sign ofloss of composite action of the beams suggesting that the shearstuds performed adequately and maintained composite actionbetween the slab and beams during the full duration of the test.

The maximum recoded deflection of the slab was 783mm,which occurred after 112min (Fig14a), which is well into thecooling stage of the fire. Fig 14b shows the time/displacementcurve for Beams 4 and 5, during the test and after one dayfollowing the test. Fig 14b also shows the deflection after onemonth once the sandbags had been removed. As shown in Fig 15,the lateral horizontal displacement of the column on gridline Dreached 2.7mm inwards into the compartment until 40min, when itthen started to move back gradually to its original position.

The deflection profile of the floor slab, coupled with thecomposite action between the beams and slab, caused rotation ofthe top flange of the steel beam. This induced a secondary

11 Recorded temperatures at mid-span of the unprotected beams 12 Recorded maximum temperatures in the unprotected beams

13(a) Recorded temperatures of the mesh above the beams

13(c) Recorded temperatures of the shear studs13(b) Recorded temperatures of the mesh betweenbeams



moment into the beam section, together with vertical shear force,leading to distortional buckling of the CBs driving the lower teelaterally out of its original plane. At this point only the upper teeprovided any strength to supporting the slab. At this stage the loadwas predominately supported by membrane action of the floorslab, corresponding to fundamental principles outlined in the Baileydesign method.

Fig 16 shows the internal beam’s residual deformation, at mid-span and towards its support, following the test. It can be seenthat the unprotected cellular beams effectively went into catenaryaction, with only the top tee contributing to supporting the load.Web post buckling, which is commonly observed in isolated small-scale fire tests, occurred around the first opening in the beamwhere the overall displacements are restricted.

Membrane action in floor slabs

The steel deck reached temperatures in excess of 900°C and wasobserved to have de-bonded from the concrete in most areas. Ata temperature of 900°C the steel deck had lost 94% of its strengthand therefore, coupled with de-bonding, did not significantlycontribute to the overall strength of the floorplate at the point ofmaximum fire severity. This corresponds to the design assumptionby Bailey where the contribution from the steel deck is ignored inthe calculation of the load capacity of the slab. However, it is worthnoting that the steel deck does have the beneficial effect ofreducing the consequence of any spalling since it ensures that anyspalled/cracked concrete stays in place, provided that the deckdoes not significantly debond and creates a large gap between thedeck and concrete. In the test a large crack occurred across theshort span of the floor slab (Figs 17a and 17b) corresponding tothe previous test observations of membrane action14, 28.

The supported concrete slab was not horizontally restrained

around its perimeter and the supporting protected perimeterbeams maintained their load carrying capacity and were subjectedto small vertical displacements. This allowed membrane action todevelop with the in-plane forces in the central region of the slabgoing into tension and in-plane equilibrium compressive forcesforming in the slab around its perimeter (Fig 17). This behaviour isanalogous to a bicycle wheel; the spokes representing tensilemembrane action, and the rim representing compressivemembrane action. For horizontally unrestrained concrete slabs afull-depth crack occurs across the shorter span, as shown in Fig17, since the strains in the reinforcement across the shorter spanare relieved as the edges of the longer side of the slab are pulledinwards as the slab deflects vertically. This behaviour has beenobserved in numerous tests and corresponds to the current designapproach to incorporate membrane action of slabs, allowing aproportion of steel beams within a floorplate to be unprotected14.

Comparison with the Bailey method

Although the slab was designed using the parametric fire designcurve outlined in BS EN1991-1-2, which provided anunconservative estimate of the fire severity, the floorplate designedusing the Bailey method supported the load for the full duration ofthe fire. The method has a number of conservative assumptions11

which compensated for the unconservative estimate of the fireseverity from the parametric fire curve.

Using the method from first principles it is possible to define afailure envelope for the slab based on the recorded temperaturesthrough the slab. As explained in previous publications11, 14 themethod is based on plastic design with change in geometry.Therefore it is not possible to follow the load–displacement curvethroughout the duration of the test but a failure envelope can bedefined. For example using the equations presented in reference 14

15 Column’s displacement recorded on the slab/beam 16 Internal beam near mid-span and end connection after fire (Beam 4)

14(a) Mid-span deflection 14(b) Deflection profile recorded on the slab/beam



we can see that if the reinforcement mesh retains its full strengththe slab will require a displacement of 771mm to enablemembrane action of the slab to support the load. Provided themesh reinforcement remained below 300°C the mesh retained itsfull strength27. At the recorded maximum temperature of the meshof 375°C the mesh has lost 5% of its strength and the requireddisplacement to support the load through membrane action needsto increase to 792mm.

The required displacement to support the load throughmembrane action of the slab alone is shown in Fig 18. Comparisonwith the recorded displacement at the centre of the slab showsthat the design method predicts the load capacity of the slab veryaccurately as the two curves merge at the maximum recordeddisplacement.

Conclusions

Full scale cellular beams were exposed to a natural compartmentfire to investigate their fire performance. The test was designedfollowing Bailey’s design method. The floorplate performedextremely well supporting the applied load for the duration of thetest and highlighted the inherent strength in the system due tomembrane action of the floor plate.

Based on the measured data it was shown that thereinforcement in the central region of the slab was under tensileforce forming an elliptical parabolic tensile mesh anchored by aconcrete compressive ring forming around the perimeter of theslab. Due to membrane action, the existence of secondary beamsto support the slab is not necessary in the fire condition and these

beams can be left unprotected.In terms of the performance of the unprotected CBs the

following conclusions can be drawn:- Due to the combined composite action of the supporting CBsand slab, distortional buckling of the CBs was the governing modeof structural failure rather than web post buckling or Vierendeelmechanism that was commonly observed in small-scale fire testson CBs in fire. - From the time when distortional buckling occurred, only the toptee of the CB’s contributed to the loading capacity of the floorplatethrough catenary action.- The CBs did not affect the membrane behaviour of the floorslab,which followed the classic behaviour as outlined in Bailey’s designmethod and supported the load for the duration of the test. - The test results compared very well with the Bailey designmethod.- The OZone program and the predicted parametric fire curve, asspecified in BS EN1991-1-2, have been shown to produceunconservative results, in terms of lower atmosphere temperaturesduring the heating phase, when compared against the test results.Further work is needed to address this unconservatisim to ensuresafe design.- The masonry wall forming the boundary of the compartmentretained its integrity despite a significant thermal gradient acrossthe wall and substantial lateral deformation. In addition, all theconnections (although protected) performed very well and showedno signs of failure.

18 Comparison between Bailey method and test results

17(a) Cracking pattern on the slab 17(b) Cracking pattern highlighting behaviour of the slab

Recorded maximum displacement at centre of the floorplate

Failure envelope from Bailey Method (Plastic design with change of geometry)



References1 Knowles, P. R.: ‘Castellated beams’, ICE Proc., 90/3, p 521–536, 19912 Dougherty, B. K.: ‘Castellated beams: a state of the art report. Technical

report’, J. South African Institution of Civil Engineers, 35/2, p12-20, 19933 Zaarour, W., Redwood, R. G.: ‘Web buckling in thin webbed castellated

beams’, J. Struct. Eng., ASCE, 8, p 860-866,19964 Vassar, O., Nadjai, A.: ‘Analytical model for the web post buckling in cellular

beams under fire’, 5th Int. Conf. Structures in fire, p 3-11, May 20085 Nadjai, A., Goodfellow, N., Ali, F.: ‘Simple calculation method of composite

cellular beams at elevated temperatures’, 5th Int. Conf. Structures in fire, p551-559, May 2008

6 Nadjai, A., Olivier, V., Ali, F.: Ahmed Allam: ‘Performance of cellularcomposite floor beams at elevated temperatures’, Fire Safety J., 42, p 489-497, Sept. 2007

7 Nadjai, A., Bailey, C. G., Ali, F.: ‘Experimental and numerical investigation oncomposite floor cellular steel beams in fire’. Steel and Composite Structures,p 673-686, Manchester, August, 2007

8 Bailey, C.: ‘Indicative fire tests to investigate the behaviour of cellular beamsprotected with intumescent coatings’. Fire Safety J., 39 (2004) p 689-709

9 Bailey, C. G., Lennon, T., and Moore, D. B.: ‘The behaviour of full-scalesteel-framed buildings subjected to compartment fires’, The StructuralEngineer, 77/8 April 1999, p 15–21

10 O’Conner, M. A, Kirby, B. R and Martin, D. M.: ‘Behaviour of a multi-storeycomposite steel framed building in fire’. The Structural Engineer, 81/2January 2003, p 27-36

11 Bailey, C. G. and Moore, D. B.: ‘The structural behaviour of steel frames withcomposite floorslabs subject to fire: Part 1: Theory’. The Structural Engineer,78/11 June 2000, p 19 – 27

12 Bailey, C. G. and Moore, D. B.: ‘The structural behaviour of steel frames withcomposite floorslabs subject to fire: Part 2: Design’. The Structural Engineer,78/11 June 2000, p 28 – 33

13 Bailey, C. G.: ‘Membrane action of unrestrained lightly reinforced concreteslabs at large displacements’, Engineering Structures, 23, p 470-483, 2001

14 Bailey, C. G., Toh, W..S.: ‘Behaviour of concrete floor slabs at ambient andelevated temperatures’. Fire Safety J., 42/ 6-7, 2007 p 425-436

15 Newman, G. M., Robinson, J. T. and Bailey, C. G.: Fire Safe Design: A NewApproach to Multi-Storey Steel-Framed Buildings (Second Edition). SCIPublication P288, The Steel Construction Institute, Ascot, 2006, ISBN 185942 169 5

16 Bailey, C. G.: ‘Steel structures supporting composite floor slabs: design forfire’. BRE Digest 462, December 2001. ISBN 1 86081 527 8

17 Advanced Structural Fire Engineering, The Institution of Structural Engineers.August 2007, ISBN 978-0-901297-46-4

18 BS 5950-1:2000: Structural use of steelwork in building. Code of practice fordesign. Rolled and welded sections. British Standards Institution, London,2000. (Withdrawn 30 March 2010, replaced by BS EN 1993-1-1:2005)

19 BS 5950-3.1:1990+A1:2010: Structural use of steelwork in building. Designin composite construction. Code of practice for design of simple andcontinuous composite beams. BSI, London, 2010. (Withdrawn 30 April2010, replaced by BS EN 1994-1-1:2004)

20 SCI P300, Composite slabs and beams using decking; Best practice fordesign and construction, 2000

21 SCI Advisory Desk, AD 325: Curtailment of transverse bar reinforcement incomposite beams with steel decking, 2008

22 NA to BS EN 1990:2002 + A1:2005; UK National Annex for Eurocode Basisof structural design. British Standards Institution, London, 2005

23 BS EN1991-1-2: 2002 Actions on structures. Actions on structures – Part1.2: General actions – Actions on structures exposed to fire. BritishStandards Institution, London, 2002

24 NA to BS EN 1991-1-2: 2002: UK National Annex to Eurocode 1: Actions onstructures – Part 1.2: General actions – Actions on structures exposed tofire, British Standards Institution, London, 2007

25 BS 7974-1: Application of fire safety engineering principles to the design ofbuildings- Part 1: Initiation and development of fire within the enclosure oforigin (Sub-system 1). Code of practice. British Standards Institution,London, 2003

26 Cadorin, J-F., Pintea, D.., Dotreppe, J-C. and Franseen, J-M.: ‘A tool todesign steel elements submitted to compartment fires – OZone V2. Part 2:Methodology and application’. Fire Safety J., 38/6, p 429-451

27 BS EN1994-1-2. Eurocode 4.Design of composite steel and concretestructures. Part 1.2. General rules. Structural fire design. British StandardsInstitution, London, 2005

28 Bailey, C. G., Toh, W. S.: ‘Small-scale concrete slab tests at ambient andelevated temperatures’. Eng. Struct., 29, 2007, p 2775-2791

Acknowledgment

The authors wish to acknowledge the supporting of the ResearchFund for Coal and Steel, Seamus O’Connor, Martin Cox andThomas Farrell from the Fire Direct Glass and Sam Kelly from thecomposite decking and Kingspan Ltd.

Registered Charity with the Charity Commission for England and Wales No.233392 and in Scotland No.SCO38263

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